Thermomechanical Actuator-Based Three-Axis Optical
Scanner for High-Speed Two-Photon Endomicroscope
Imaging
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Citation
Chen, Shih-Chi, Heejin Choi, Peter T. C. So, and Martin L.
Culpepper. “Thermomechanical Actuator-Based Three-Axis
Optical Scanner for High-Speed Two-Photon Endomicroscope
Imaging.” Journal of Microelectromechanical Systems 23, no. 3
(June 2014): 570–578.
As Published
http://dx.doi.org/10.1109/jmems.2013.2287708
Publisher
Institute of Electrical and Electronics Engineers (IEEE)
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Author's final manuscript
Accessed
Thu May 26 12:40:06 EDT 2016
Citable Link
http://hdl.handle.net/1721.1/97404
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http://creativecommons.org/licenses/by-nc-sa/4.0/
NIH Public Access
Author Manuscript
J Microelectromech Syst. Author manuscript; available in PMC 2015 February 09.
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Published in final edited form as:
J Microelectromech Syst. 2014 May 29; 23(3): 570–578. doi:10.1109/JMEMS.2013.2287708.
Thermomechanical Actuator-Based Three-Axis Optical Scanner
for High-Speed Two-Photon Endomicroscope Imaging
Shih-Chi Chen,
Department of Mechanical and Automation Engineering, The Chinese University of Hong Kong,
Hong Kong
Heejin Choi,
Department of Mechanical Engineering, Massachusetts Institute of Technology, Cambridge, MA
02139 USA
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Peter T. C. So, and
Department of Mechanical Engineering, and the Department of Biological Engineering
Massachusetts Institute of Technology, Cambridge, MA 02139 USA
Martin L. Culpepper
Department of Mechanical Engineering, Massachusetts Institute of Technology, Cambridge, MA
02139 USA
Shih-Chi Chen: scchen@mae.cuhk.edu.hk; Heejin Choi: choihj@mit.edu; Peter T. C. So: ptso@mit.edu; Martin L.
Culpepper: culpepper@mit.edu
Abstract
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This paper presents the design and characterization of a three-axis thermomechanical actuatorbased endoscopic scanner for obtaining ex vivo two-photon images. The scanner consisted of two
sub-systems: 1) an optical system (prism, gradient index lens, and optical fiber) that was used to
deliver and collect light during imaging and 2) a small-scale silicon electromechanical scanner that
could raster scan the focal point of the optics through a specimen. The scanner can be housed
within a 7 mm Ø endoscope port and can scan at the speed of 3 kHz × 100 Hz × 30 Hz along three
axes throughout a 125 × 125 × 100 μm3 volume. The high-speed thermomechanical actuation was
achieved through the use of geometric contouring, pulsing technique, and mechanical frequency
multiplication (MFM), where MFM is a new method for increasing the device cycling speed by
pairing actuators of unequal forward and returning stroke speeds. Sample cross-sectional images
of 15-μm fluorescent beads are presented to demonstrate the resolution and optical crosssectioning capability of the two-photon imaging system.
Index Terms
Thermomechanical actuators; mechanical frequency multiplier; two-photon excitation;
endomicroscope; optical scanner
© 2013 IEEE.
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I. Introduction
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The development of a two-photon excitation (TPE) endomicroscope enables real-time subcellular resolution volumetric imaging with an imaging depth of 200–500 microns. This
instrument has the potential to serve as a powerful tool for clinicians and pathologists to
improve their diagnostic accuracy and efficiency and ultimately realize the concept of
“optical biopsy”.
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Typical TPE microscope raster scans a point focus in 3D to produce a volumetric image.
The design of a TPE endomicroscope requires miniaturization fitting a three-axis scanner
within the envelope of an endoscope. The scanners must operate at relatively high speed,
e.g. 1–3 kHz to achieve a clinically acceptable frame rate. The scanning range of each
individual micro-scanner defines the field of view of the endomicroscope that is typically on
the order of tens to hundreds of microns on a side. As such, selections of micro-actuators/
scanners that are compatible with silicon fabrication process and simultaneously fulfill the
range and speed requirement are critical. Existing confocal or TPE endomicroscope systems
utilize piezoelectric actuators [1]–[4] or electrostatic actuators, e.g. MEMS mirrors [5]–[8]
for actuation and scanning. These actuators operate at high voltage, and thus present a safety
concern for in vivo operations. A few thermomechanical actuator-based MEMS scanners
were used in endoscopic optical coherence tomography imaging system with limited
scanning speed, i.e. less than 5 frames per second [9], [10]. A low voltage and low power
high-speed scanning system has yet to be developed.
In this work, we present the modeling, design, fabrication, and characterization of a highspeed three-axis optical scanning system for TPE endoscopy application—all with
thermomechanical actuators (TMA). TMAs are known for their high force/power density
but with limited bandwidth; we show by applying the “geometric contouring” and the
“mechanical frequency multiplication (MFM)” design concept, a TMA-based high-speed
endoscopic scanning system may be developed to generate real-time two-photon volumetric
images with low operation voltage, i.e. less than 5V and low power level, i.e. 150mW.
Although a TMA runs at high temperature, e.g. 1000 K, it causes no danger with proper
insulation since the heated volume is on the micron scale and this energy dissipates quickly
as long as the power level is low.
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1) Functional Requirements—A TPE endomicroscope generates tissue images by
scanning the focused point of light throughout a volume. For in vivo imaging, the scanning
process should be fast such that the physiological motions, such as heart beat, of the subject
do not compromise the fidelity of the image. Depending upon the number and the size of
optical sections, this constraint will set the scanning speed and range. To satisfy the
aforementioned requirements and constraints, the endoscopic scanner should operate at 3
kHz, 30 Hz, and 2 Hz for the X, Y, and Z (optical) axis respectively. With a 2D frame rate
of 30Hz and a 3D image rate of 2Hz, motion artifacts will cause variances in the axial
spacing between 2D frames but each 2D frame will remain mostly motion artifact free. This
inaccuracy in the axial direction is often tolerable because histopathologists base their
analyses primarily on 2D image morphological features. The focused laser should scan 100
microns in all three axes to sample a sufficiently large tissue volume. The envelope of the
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scanner should reside within a 7mm Ø endoscope port. These clinically driven functional
requirements are summarized in Table I. The preliminary ex vivo experiments presented
later in this paper were scanned at 1kHz, 2Hz, and 0.1Hz in each axis in order to increase the
signal to noise ratio. Note for some in vivo experiments, where weak endogenous
fluorescent signals are used to generate images, the frame rate is actually limited by the
signal strength rather than the mechanical scan speed.
II. Design for the Distal Optics of a TPE Endomicroscope
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Figure 1 shows the optical design of the TPE endomicroscope system. In this system,
ultrafast optical pulses from a Ti:Sappire laser with 100fs pulse width, 80MHz repetition
rate at the center wavelength of 780nm (MaiTai, Spectra-Physics, Mountain View, CA)
were used as the excitation light source. A double clad photonic crystal fiber (DCPCF)
(DC-165-16-Passive, Crystal Fibre A/S) was used to deliver both the excitation and
emission light. (DCPCF allows a single mode excitation beam delivery through its core and
emission beam delivery through the core and inner clad, thereby simplifying the
endomicroscope design and increasing the detection efficiency.) A pre-chirping unit was
used to precompensate the linear pulse dispersion induced by DCPCF and maximize the
two-photon excitation efficiency, which is inversely proportional to the pulse width of the
excitation beam. After propagating through the DCPCF, the pulsed excitation beam was
guided to the silicon scanning bench, going through a prism and lastly a gradient index
(GRIN) lens as the objective lens (GRINTECH: GT-LFRL-100-017; NA = 0.5). The
emission from the focal point in the sample then traveled back through the optical train, and
was guided to a photo-multiplier tube (PMT) detector via a dichroic mirror (DM). The
relative positions among different optical components were optimized by ray-tracing
analysis with Zemax (Radiant Zemax, Redmond, WA) in order to maximize resolution and
minimize aberrations.
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To obtain volumetric images, the DCPCF, prism, and GRIN lens need to be scanned in the
X, θX, and Z direction respectively. The DCPCF will generate the X-axis scanning motion
by a TMA-based fiber resonator; the prism and the objective GRIN lens will generate the Yaxis and Z-axis scanning respectively by TMA-actuated shuttles on the silicon scanning
bench. Note that the Y-axis scan is performed via oscillating the prism about the θX axis; the
100 micron requirement is equivalent to a 2° angular scanning motion.
III. TMA-Based Two-Axis Silicon Scanning Bench
This section presents the modeling, design, and experimental characterization of the twoaxis endoscopic scanner. Figure 2A shows the optical system on the silicon optical bench.
The scanner is comprised of a GRIN lens shuttle and a prism shuttle, each integrated with
guiding flexures and TMAs. The GRIN lens and the prism are of millimeter scale; the
optical bench provides micrometer-level precision alignment for the lenses. The
performance of the chevron TMAs are optimized through the geometric contouring
technique [11]. High-speed operation may be achieved through the application of the highspeed pulsing technique [12].
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A. Flexural Bearing for Motion Guidance and Amplification
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The scanner consists of two sets of flexural bearings with chevron TMA trains, i.e. the
GRIN lens shuttle and the prism shuttle. The chevron TMAs were selected over parallel
TMAs because for a given footprint they were able to run in parallel and generate larger
force output without sacrificing the stroke. TMAs used in the system were optimized via
“geometric contouring” technique with the following objective: (1) minimize power
consumption and (2) maximize stroke. This was desired because the TMA trains on the
scanner operate at a moderate speed of 30Hz and 2Hz, but require a rather large stroke of
100 microns. The speed is not critical here as with the pulsing technique the TMA train is
able to achieve a 10-times faster cycling speed [12]. As the flexure mechanism may easily
provide more than 10X motion amplification, the required stroke for individual TMA was
set to be 10 micron. The optimal design parameters were selected by using the contoured
chevron TMA performance charts provided in [11]. The TMA performance for the GRIN
lens shuttle and the prism shuttle was modeled and is summarized in Table II.
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1) Translation Flexural Bearing Concept—The flexural bearing that was actuated by
the TMA trains provides precision guidance and reduced parasitic motion for active optical
components. As shown in Figure 2A, the flexural bearing that carries the objective GRIN
lens generates linear motion and travels 100 microns. This flexural bearing consists of a
two–stage chevron amplification mechanism that was driven by two sets of chevron TMA
trains. The symmetric design is used to prevent lateral parasitic motion. Note that only the
chevron TMA train was used as an actuator. The interlinking chevron flexure was for
transmitting/amplifying mechanical motion, and did not experience internal heat generation.
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2) Rotary Flexural Bearing Concept—The second flexural bearing carried the prism
and generated rotary motion about the instant center over +2.50°. This rotary bearing was
driven by the two-stage chevron amplification mechanism and chevron TMA train, shown in
Figure 2A. The shuttle that carries the prism was part of a flexural four-bar mechanism. The
instant center of this four-bar mechanism was designed to be on the reflecting point of the
prism’s inner surface. Parasitic translational motion should therefore be minimized during
the rotation. By supplying the TMA train a constant power and varying the angle between
the two supporting flexures of the four-bar mechanism, the optimal angle of the four-bar
mechanism can be found. As shown in Figure 3, at 70° the flexural rotary bearing yields 6
nm of parasitic motion at full stroke (2.5°). As the GRIN lens and prism weigh 1.5 and 1.2
grams respectively, the gravity effects on both shuttles were modeled and proven to be
below a level that would cause practical concern.
B. Mechanical Impedance Matching
With the optimized flexural four-bar mechanism, we now model the cascaded chevron
flexure in order to optimize its transmission ratio, defined as the ratio of the output
displacement to the input displacement.
The transmission ratio, axial stiffness, and lateral stiffness of a chevron flexure, are all
functions of the angle (θ) between the chevron beam and the horizontal line shown in Figure
4A. In Figure 4B, the transmission ratio and the lateral stiffness of a chevron mechanism
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drop sharply, while the axial stiffness increases gradually, as the angle (θ) increases. It is not
uncommon for designers to attempt to cascade two chevron mechanisms at small angles in
order to receive a large transmission ratio. In fact, this design will result in a low
transmission ratio because of the mismatch between the axial and lateral stiffness of the first
(TMA) and second (amplification flexure) chevron mechanisms that are shown in Figure 5.
A basic concept that may be used for optimizing the energy transfer from the first chevron
TMA to the second chevron mechanism is to match their mechanical impedance, i.e. design
the chevron flexures so that the axial stiffness (KA1) and lateral stiffness (KL2) of the first
and second chevrons are equal. There are two ways to adjust the relative values of KA1 and
KL2 for mechanism optimization: (1) adjust angles of the first and second TMAs, i.e. θ1 and
θ2, as shown in Figure 4B and Figure 5, or (2) use many chevron TMAs in parallel, where
the effect of increased axial stiffness helps to increase the transmission ratio as shown in
Figure 4C and Figure 5, where N represents the number of parallel TMAs. Although a large
transmission ratio may be obtained through the second option, this may result in a large
power requirement to run many TMAs simultaneously.
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As shown in Figure 5, the cascaded chevron mechanism may be modeled as a two-spring
system, where spring constants KA1(θ1) and KL2(θ2) represent the axial stiffness and the
lateral stiffness of the first and second chevron mechanism respectively. The variable XIN
represents the displacement of a chevron TMA (first chevron mechanism), and X is the
output displacement of the cascaded system. The relationship between X and XIN is shown in
Equation (1). The overall transmission ratio of the cascaded mechanism is shown in
Equation (2). The stroke of cascaded system is the product of the input displacement and the
transmission ratio, as shown in Equation (3).
(1)
(2)
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(3)
Equation (2) was used to set a transmission ratio as a function of θ1 and θ2. Figure 6A
presents the optimal transmission ratios as a function of the number of parallel TMAs. The
normalized mechanism stroke, i.e., power efficiency, as a function of number of TMAs, may
be obtained by dividing the optimal transmission ratio by the number of TMAs. Figure 6A
indicates that more parallel TMAs lead to a larger transmission ratio. Unfortunately, there is
a corresponding decrease in power efficiency, as shown in Figure 6B. A minimum number
of TMAs should be used if the optimization objective is to minimize the power
consumption.
A general approach that may be used to obtain an efficient design is summarized below:
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Step 1: Determine the required stroke (S) from functional requirements.
Step 2: Determine the required displacement (d) of an individual TMA.
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Step 3: Determine the required optimal transmission ratio (TR) by dividing the system
stroke by the individual TMA’s displacement (TR = S/d).
Step 4: Obtain the corresponding values for θ1 and θ2 for the optimal transmission ratio
from the transmission equation.
Note that using Equation (2), we can obtain surface plots of the transmission ratio as a
function of θ1 and θ2 with different number of TMAs, shown in Figure 7A. When the
required transmission ratio is determined, optimal values of θ1 and θ2 can be found through
the surface plot, as shown in Figure 7B.
C. Modeling and Optimization
In this section, we discuss how best design is achieved. We first consider two optimization
objectives: (1) temperature minimization design and (2) power minimization design. The
GRIN lens shuttle is used as an example to demonstrate the design process.
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As the scanner is designed for in vivo operation, one may desire a reduced operating
temperature. The design process is: (1) determine the TMA’s maximum operating
temperature from functional requirements, e.g., 150°C, (2) optimize the TMA’s
performance, e.g. stroke, at the specified temperature by the geometric contouring method,
and (3) obtain the required number of TMAs and the relative optimal transmission ratio
from Equation (2). This approach leads us to a 20-TMA design for 100 micron output
displacement at 150°C, where the transmission ratio is 45.1, θ1 = 2.0°, and θ2 = 2.0°.
For the power minimization, we can set the maximum operating temperature to 1200 K as a
TMA operates more efficiently at high temperature. Following the same design process, we
find with the power minimization approach, we can achieve 100 micron output displacement
with merely 2 parallel TMAs at 125 mW power consumption. The related transmission ratio
is 28.1, for θ1 = 2.5°, and θ2 = 3.6°.
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Although to minimize operation temperature or power may seem reasonable, each of them
has its drawbacks. For the low temperature design, the combined power consumption of the
GRIN lens and prism shuttle on a single chip may be too high, i.e. 1.2W, making this device
inefficient. For the low power design, the maximum stroke at 100 micron will be reached at
1200°K which leaves the device no room for additional displacement. Based on these
analyses, a better approach is to set a fixed number of TMAs and optimize transmission
ratio. Additionally, the use of more parallel TMAs increases the shuttle’s structural integrity.
For the best design, we choose the number of TMAs (N) to be 5. Accordingly, this design
easily satisfies the stroke requirement, achieving a 100 micron output displacement at 160
mW power consumption, and a maximum stroke of 400 micron at 1200°K. The transmission
ratio is 43.8, θ1 = 2.2°, and θ2 = 2.8°. The best design was thus chosen for fabrication and
testing.
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The design parameters and simulated performance of the GRIN lens shuttle and the prism
shuttle for three different design objectives are summarized in Table III.
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IV. Experimental Results and Discussion
Figure 8 shows frequency spectrums of the GRIN lens and prism shuttle (note experimental
data in this section were obtained with lenses loaded). The simulated first and second mode
of the GRIN lens shuttle generate in-plane and out-of-plane motion at 1,527 Hz and 7,121
Hz respectively. The measured in-plane resonant peak is at 1,508 Hz, which matches the
simulated value within 1.3%. For the prism shuttle, the simulated values of first and second
mode are 3,205Hz and 7,096Hz respectively; the measured corresponding resonant peaks
(3,124 Hz and 7320 Hz) match the simulated values with errors of +2.6% and −3.0%
respectively, as shown in Figure 8B. The accuracy of the frequency modeling is attributed to
the accurate geometry values that were obtained via measurement within the SEM. The
frequency results also indirectly indicate that there were no hidden cracks or broken flexures
in the device.
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Figure 9 shows the step response of the GRIN lens and prism shuttle. The measured rise
time and fall time for the GRIN lens shuttle were 0.46 and 0.38 second respectively. These
values show an error of −6.5%. For the prism shuttle, it was found that the rise time and fall
time were 0.13 and 0.09 second respectively. The worst case error between these values and
the predicted values was +7.7%. We also found the cycling speed of the GRIN lens shuttle
fulfills the requirements, i.e. to scan at 2 Hz, before the pulsing technique was applied.
Details of the TMA modeling approach may be found in [11], [12].
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Figure 10 shows the results for static displacement of the GRIN lens and prism shuttle,
demonstrating they fulfill the stroke requirement. Measured positions were plotted as a
function of command position in Figure 10A and Figure 10B. In both experiments, the
measured displacements show a gradual divergence from commanded values at elevated
temperature, where a +6.7% error was observed at full stroke for the GRIN lens shuttle. This
divergence was due in part to the temperature dependence of electrical resistivity and the
differences in the manner in which electrical properties of dopants (phosphorous and arsine)
change with temperature. Assuming the thermal conductivity and coefficient of thermal
expansion (CTE) of our arsine-doped wafer were accurate, the resistivity would have a 7%
error at elevated temperatures. Note that the fabrication errors were not included in the error
analysis as all critical dimensions of the devices were measured in the SEM after they were
fabricated.
The match between the simulated and experimental results in both dynamic and static
measurements gives confidence that our parametric models may be used by other designers
to predict device performance with less than 7% error.
V. Mechanical Frequency Multiplication for Fiber Resonation
This section presents the modeling, design, and experimental characterization of the TMAbased fiber resonator based on the concept of “mechanical frequency multiplication” [13], as
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shown in Figure 11. The MFM resonator excites the DCPCF at its resonant frequency, i.e.
1–3kHz, in order to generate a linear scanning pattern for the high-speed X-axis.
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An MFM device consists of three components: (1) a pulse-generating TMA pair, (2) a
decoupling flexure, and (3) a main stage that is connected and guided by flexures. The
pulse-generating pair utilizes the fact that TMAs possess different forward and return stroke
speeds. The concept of MFM is to use the high–speed portion of the one actuator to rapidly
achieve half–cycle motion and then use the opposing actuator to rapidly return the last half
cycle. If N actuator pairs are placed in parallel, time delayed signals may be used to drive
each set with a delay, thereby increasing the cycling frequency by N.
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Figure 11 presents schematics of the MFM concept and its operation principle. Figure 11A
shows examples of a pulse generation TMA pair, where a high frequency mechanical pulse
output may be generated by pairing two TMAs, “α” and “β”, in opposing directions and
supplying them with time-delayed power input α and β respectively; Figure 11B shows short
pulses generated by a TMA pair—although TMA α and β cool down slowly, the resultant
motion for the TMA pair forms a pulse. Figure 11C and 11D shows the temperature and
displacement effects of combining multiple pulse pairs, achieving high cycling frequency. It
is worth to note that the mechanical pulse width is not limited by cooling time/process. It is
only limited by the dynamic characteristics of the actuator pair that is the pair’s resonant
frequency. An actuator system with high bandwidth may be constructed if many of the TMA
pairs act in parallel to drive a common stage with time delays in their pulses.
A. MFM Fiber Resonator Design
The design concept of the MFM fiber resonator is shown in Figure 12. Two pulse-generation
chevron TMA pairs are located at either side of the main stage, which is the most basic form
of an MFM system. In this design, four TMAs also function as (1) the motion guiding
flexures and (2) the coupling flexures that transmit their motions to the stage. This flexure
concept provides the MFM with a high mechanical resonance frequency (17.7 kHz). The
device was designed to fit within a 2 × 2 mm2 envelope. The fourTMA design was able to
achieve 4 times the cycling frequency of one of its constituent TMAs.
B. TMA Selection and Design
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Geometric contouring designs were applied to the TMAs in the MFM fiber resonator to
enhance stroke, efficiency and power consumption. It is known that the forward and return
speed ratio of a contoured TMA may be controlled by either the input command or the
design parameters for a contoured beam [12]. It is therefore important to design each
constituent contoured TMA of the four-TMA MFM system so that its fall time (cooling
time) is equal to, or larger than, four times its rise time (heating time). The MFM system
will perform more efficiently when this requirement is met. Accordingly, the contoured
TMAs of the MFM system were then designed and optimized based on this objective and
the static/dynamic TMA performance charts provided in [11], [12]. The finalized design
parameters of TMAs are listed in Table IV, where LS, LL, w′ and we are defined and
discussed in detail in [11], [12].
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C. Fiber Resonance Experiment
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Figure 13 shows the results of the resonance on the fiber’s tip. The images were acquired via
the CCD camera. Figure 13A shows a still image of the fiber’s tip before the MFM was
energized, and Figure 13B shows the fiber movement patterns that were generated by
actuating the MFM with properly coordinated actuation signals. The amplitude of the
scanning range was estimated to be 125±2 micron, which satisfies the functional
requirements. The accuracy of the amplitude was ascertained by pixel-counting the image
obtained from the CCD camera, where each pixel in the image equals to 0.5 μm.
VI. Microfabrication Processes
The microfabrication process, shown in Figure 14, for both the silicon optical bench and the
MFM fiber resonator are described below:
Step 1: The process starts with a SOI wafer (device layer: 200 micron; resistivity =
0.001 ohm-cm).
Step 2: Deposit and pattern 300 nm aluminum contacts for the device through
sputtering.
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Step 3: Pattern the device structure with deep reactive ion etching (DRIE).
Step 4: Target–mount the device wafer onto another silicon wafer via photoresist.
Step 5: Release the device with a backside through–etch using DRIE.
Step 6: Release the mounted device wafer and clean the photoresist/residues.
VII. Imaging Experiments and Preliminary Results
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Custom-developed data acquisition software and control electronics were constructed to
synchronize the different scanning axes with the photon counting circuitry monitoring the
optical signal from the photomultiplier tube. 3D images were reconstructed computationally
by correlating the measured optical signal strength with the known scanner positions.
Characterization of individual scanner was first performed to confirm the endomicroscope’s
field of view as well as the reliability of the system, as discussed in previous sections. To
demonstrate 3D optically sectioned imaging capability of the TPE endomicroscope system
and to characterize the resolution of the system, fluorescent beads of different sizes,
including 5, 10, 15 microns, were used for preliminary imaging experiments. The results
indicate that the endomicroscope has a lateral resolution of approximately 1.0 micron and
axial resolution of 8.0 micron with a penetration depth of 100 micron (limited by the range
of Z scan). Figure 15 shows an example of fluorescent bead cross section images obtained
through the two-photon endomicroscope. The imaging results confirm the effectiveness of
TMA-based micro-scanning system.
VIII. Conclusion
We have presented the design, modeling, fabrication and experimental characterization of a
silicon optical scanner and a MFM fiber resonator—all based on TMAs. The design method
of “geometric contouring” was used in optimizing TMAs in all devices. Experimental results
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show the fabricated optical scanner, containing the GRIN lens scanner and the prism
scanner, achieved 100 micron and 2.5° scanning range and the MFM resonator achieved 125
micron scan range, satisfying all functional requirements at a power level of approximately
150mW and a 5V operation voltage; this presents a minimized safety risk for in vivo
experiments. Our parametric models also match well with the experimental results both in
static and dynamic measurements, providing designers a fast and accurate design approach
rather than entirely resorting to time-consuming finite element analysis. (Note FEM was still
used for calculating the natural frequencies and system mode shapes as well as for finetuning the final design parameters.) These results enable designers to systematically obtain
the optimal mechanical design for their applications. In the end, we present the preliminary
imaging results showing TPE optical cross-sectional images may be reliably obtained via the
TMA-based scanners. The endomicroscope is characterized to have an in-plane resolution of
1.0 micron and axial resolution of 8.0 micron. The results show great potential of using the
TPE endomicroscope for real-time disease diagnosis. Future efforts will be focused on
applying the TPE endomicroscope for in vivo animal studies.
Acknowledgments
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This work was supported in part by the National Institutes of Health under Grant 1-R21-CA118400-01: Design of a
Non-Linear Endomicroscope Biopsy Probe, and the Chinese University of Hong Kong Direct Grant 2050495:
Design of a Two-Axis High-Speed Scanning Mirror for Advanced Non-Linear Microscopy. The work of H. Choi
and P. T. C. So was supported in part by the National Institutes of Health under Grant 9P41EB015871-26A1, Grant
5R01EY017656-02, Grant 5R01 NS051320, and Grant 4R44EB012415-02, in part by the National Science
Foundation under Grant CBET-0939511, in part by the Singapore-MIT Alliance 2, in part by the MIT SkolTech
initiative, in part by Hamamatsu Corporation, and in part by the Koch Institute for Integrative Cancer Research
Bridge Project Initiative.
We would also like to thank Prof. Dennis Freeman (MIT EECS) for granting us free access to his Computer
Microvision System.
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Biographies
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Shih-Chi Chen received his B.S. degree in Mechanical Engineering from the National
Tsing Hua University, Taiwan, in 1999. He received his S.M. and Ph.D. degrees in
Mechanical Engineering from the Massachusetts Institute of Technology, Cambridge, in
2003 and 2007, respectively. Following his graduate work, he entered a post-doctoral
fellowship in the Wellman Center for Photomedicine, Harvard Medical School, where his
research focused on biomedical optics and endomicroscopy. He is currently an Assistant
Professor in the Department of Mechanical and Automation Engineering at The Chinese
University of Hong Kong (CUHK). His research interests include precision engineering,
biomedical optics, microsystem design, and nanomanufacturing. Prof. Chen is a Member of
the American Society of Mechanical Engineers (ASME) and the American Society for
Precision Engineering (ASPE). He is the recipient of a 2003 R&D 100 Award for the design
of a microscale six-axis nanopositioner.
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Heejin Choi received his B.S. degree in Mechanical Engineering from Pohang University of
Science and Technology, South Korea, in 1999. He received his M.S. degree in Mechanical
Engineering from Korea Advanced Institute of Science and Technology, South Korea, in
2003. He is currently a Ph.D. degree candidate in Mechanical Engineering at Massachusetts
Institute of Technology, Cambridge, MA, USA. He was the recipient of the best poster
presentation award at the 2008 Optical Society of America (OSA) biomedical optics topical
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meeting. His research is focused on the optical imaging and spectroscopy instrumentation
for the medical and biological applications.
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Peter T. C. So is a professor in the Department of Mechanical and Biological Engineering
in the Massachusetts Institute of Technology. Prior to joining MIT, Peter So obtained his
B.S. from Harvey Mudd College in 1982 and Ph.D. from Princeton University in 1992. He
subsequently worked as a postdoctoral associate in the Laboratory for Fluorescence
Dynamics in the University of Illinois in Urban-Champaign. His research focuses on
developing high resolution and high information content microscopic imaging instruments.
These instruments are applied in biomedical studies such as the noninvasive optical biopsy
of cancer, the mechanotransduction processes in cardiovascular diseases, and the effects of
neuronal remodeling on memory plasticity. Peter So is currently the Director of the MIT
Laser Biomedical Research Center, an NIH NIBIB P41 research resource and the Program
Chair of the Computational System Biology Program within the Singapore-MIT Alliance 2.
Martin L. Culpepper received his B.S. (1995) in mechanical engineering from Iowa State
University, Ames, IA; and his M.S. (1997) and Ph.D. in mechanical engineering (2000)
from the Massachusetts Institute of Technology, Cambridge, MA.
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He worked as a consultant at Teradyne from 2000–2001 before joining the faculty of
mechanical engineering at the Massachusetts Institute of Technology, Cambridge, MA. He
is currently the Director of the Precision Compliant Systems Laboratory and the Associate
Director of the MIT Laboratory for Manufacturing and Productivity.
Prof. Culpepper is a Fellow of the American Society of Mechanical Engineers (ASME) and
a member of the American Society of Precision Engineers (ASPE). He was the recipient of a
2004 NSF PECASE Award (Nanomanufacturing), two R&D 100 awards, and a TR100
award. His research is focused on the design of meso/micro/nano-scale equipment and
instruments for nanomanufacturing.
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Fig. 1.
Optical configuration of the TPE endomicroscope system.
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Fig. 2.
Schematics (A) and fabricated (B) 2-axis endoscopic scanner with integrated GRIN lens,
prism and TMA actuators.
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Fig. 3.
Flexural rotary bearing optimization. Note that at 70° the ratio of parasitic displacement over
actuator displacement is 2.9 × 10−3 [mm/mm].
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Fig. 4.
Transmission ratio, axial, lateral stiffness of a chevron mechanism as a function of angle and
the number of chevron TMAs. A: Contoured chevron TMA model, where θ is the TMA
angle, KA and KL are the axial and latteral stiffness of the TMA; B: Characteristics of a
single chevron mechanism; C: Increased axial stiffness by devising 10 or 20 chevron TMAs
in parallel.
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Fig. 5.
Two–spring model for the cascaded chevron mechanism.
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Fig. 6.
A: Optimal transmission ratio as a function of number of TMAs; B: Power efficiency as a
function of number of TMAs.
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Fig. 7.
Transmission ratio surface plot as a function of θ1, θ2, and number of TMAs (N).
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Fig. 8.
Measured frequency spectrum and for GRIN lens shuttle and prism shuttle.
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Fig. 9.
Step response of GRIN lens shuttle and prism shuttle.
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Fig. 10.
Displacement vs. input command plot for GRIN lens shuttle and prism shuttle.
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Fig. 11.
Schematics of the mechanical frequency multiplication concept and its operation principle.
A: Examples of pulse generator pairs. B: Short pulses generated by pulse pairs. C: Max.
temperature on individual pulse pair. D: Multiple pulse pairs are combined to form an MFM.
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Fig. 12.
Schematics and image of a MFM fiber resonator with fiber mounted. A: Schematics and
image. B: MFM device in operation.
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Fig. 13.
CCD images of MFM resonator in operation at 965 Hz, generating a line scan of 125
micron. A: Still image of the fiber before actuation. B: Fiber resonated by the MFM system.
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Fig. 14.
Microfabrication process for silicon optical bench and MFM fiber resonator.
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Fig. 15.
A stack of fluorescent beads images with a field of view of ~ 50 × 60 microns. Each bead is
of 15 micron in diameter and each frame is 3 micron apart in Z-axis.
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TABLE I
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Functional Requirements and Constraints of the Endoscopic Scanner
X-axis
Y-axis (θx)
Z-axis
100 μm
100 μm (2°)
100 μm
Speed (in vivo)
3 kHz
30 Hz
2 Hz
Speed (ex vivo)
1 kHz
2 Hz
0.1 Hz
Range
Envelope
Within a 7mm diameter tube
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TABLE II
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Simulated Performance of Contoured TMA for the Optical Scanner
TMA for th GRIN lens shuttle
Stroke
10 μm
Operating voltage
2.8 V
Operating temperature
580 K
TMA bandwidth
8 Hz (before pulsing)
Max. force output
10 mN
TMA for the prism shuttle
Stroke
14 μm
Operating voltage
2.8 V
Operating temperature
580 K
TMA bandwidth
16 Hz (before pulsing)
Max. force output
8 mN
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TABLE III
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Summary of Three Scanner Designs for Different Objectives
Design I
Design II
Design III
Low temperature
Low power
Best design
Transmission for GRIN shuttle
45.1
23.8
43.8
θ1
2.0°
2.5°
2.8°
θ2
2.0°
3.6°
2.2°
Stroke
100 μm
100 μm
100 μm
Power (at full stroke)
700 mW
125 mW
160 mW
Transmission for prism shuttle
42.5
11.5
18.8
θ1
2.0°
5.0°
4.5°
θ2
2.5°
5.5°
4.5°
Stroke
2.0°
4.0°
8.0°
500 mW
50 mW
125 mW
Parameters
Power (at full stroke)
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TABLE IV
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Design Parameters of Contoured TMAs in MFM Resonator
Contoured TMA design parameters for MFM fiber resonator
LS/LL
1/8
we
9.25 μm
L/2LL
5/4
L
1000 μm
w′
1/2
b
200 μm
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